# Challenges in Bioenergy Production from Sugarcane Mills in Developing Countries: A Case Study

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## Abstract

**:**

## 1. Introduction

Country | Power mode | Configuration | Use of trash | Surplus electricity (kWh/ton of cane) |
---|---|---|---|---|

Brazil | BPST | 22 bar, 300 °C | No | 0–10 |

Brazil | BPST | 42 bar, 440 °C | No | 20 |

Brazil | BPST | 67 bar, 480 °C | No | 40–60 |

Brazil | CEST | 65 bar, 480 °C | Yes (50%) | 139.7 |

Brazil | CEST | 105 bar, 525 °C | Yes (50%) | 158 |

India | CEST | 67 bar, 495 °C | No | 90–120 |

India | CEST | 87 bar, 515 °C | No | 130–140 |

## 2. Methods

#### 2.1. Case Study: A Colombian Sugarcane Milling Plant

**Figure 1.**Process in a sugarcane milling plant [10].

#### 2.2. Electrical Energy and Steam Requirements in the Sugar and Bioethanol Plant Production

^{2}= 0.756 which, according to the ISO 500001:2011 [13], is a sufficient indicator to establish the relationship. The study based on the normative outlined that 71.5% of the plant electrical energy demand was not related to the production [13]. The study showed that for the alcohol plant there was no relationship between milling rate and electrical consumption. For this reason, the electrical consumption of the PAC was a constant equal to 50,360 kWh/day. Total electrical energy consumption of the plant is the sum of the mill rate-depending part and the constant one (PAC). The electrical power requirement (P

_{el}) is then:

_{vap}), which is translated into heat power need Q

_{need}of the whole plant with Equation 4:

_{vap}= 20.32 × mr

^{−0.453}×1000 / (3600 × 2.205) (kg/s)

_{need}= 234.86 × mr + 54546 (kW)

^{3}/day of ethanol.

#### 2.3. Cogeneration

_{mill}, P

_{shred}) were assumed as constant (P

_{mill}= 12.7 (kW/tfh-mill) P

_{shred}= 34.98 (kw/tfh)). The steam mass rates of turbo-generators, the mills’ turbines, the shredders and crushers are calculated based on the mass and energy balances. The turbines connected to turbo-pumps and the turbo-fans are related to the steam flowing to the boilers and are modeled as constant at their nominal power. With these assumptions the two blocks of the scheme that represent the turbines can be modeled.

_{output}), the blade to shaft efficiencies η

_{BS}, and the thermodynamic points at which they work (h

_{in}, h

_{out}), steam mass rates (m

_{vap}) are calculated with Equation 5:

_{H2.7}, total steam mass rate at turbines outlet m

_{VAP-T}, specific enthalpy of total steam mass rate at turbines’ outlet h

_{out-TURB}, steam mass rate crossing the 21.7 to 2.7 bar reducers m

_{VAP-R21}, specific enthalpy of steam in the head of 21.7 bar h

_{H21}, specific enthalpy of 2.7 bar saturated steam h

_{SAT 2.7bar}, and the specific enthalpy of tempering water h

_{W}, the following expressions were used to calculate the required mass rate of tempering water m

_{TW}:

_{VAP-R21}) bar is different from zero if the steam requirement of the head of 2.7 bar m

_{VAP-N2.7}is greater than the steam flowing through the blocks of the turbines (m

_{VAP-TURB}):

_{VAP-R21}= m

_{VAP-N2.7}– m

_{VAP-TURB}– m

_{TW}(kg/s)

_{VAP-R28}) is almost constant at different milling rates because, due to its higher efficiency, Boiler C is kept at full capacity (cap

_{Bc}) and the steam requirement of the head of 28.6 bar is almost constant too. Then, the steam flowing to this reduction station is given by the subtraction of this head requirement (m

_{VAP-N28}) from the Boiler C capacity (Equation 9). The production on Boilers A and B (m

_{VAP-B-ab}) then covers the other part of the 21.7 bar demand (m

_{VAP-N21}) which is given by the requirement of the 21.7 bar turbines (m

_{VAP-TURB21}), the steam flowing through the reduction station, and by the PAC’s requirement (m

_{VAP-PAC}). Equations 9 to 11 provide the steam mass balance at pressure reducer at Boilers A and B, while with Equation 12 the energy balance on the 21.7 bar head is computed:

_{VAP-R28}= cap

_{Bc}– m

_{VAP-N28}(kg/s)

_{VAP-N21}= m

_{VAP-TURB21}+ m

_{VAP-R28}+ m

_{VAP-PAC}(kg/s)

_{VAP-B-ab}= m

_{VAP-N21}– m

_{VAP-R28}(kg/s)

- Temperature and pressure of Boiler C: T
_{Bc}= 410 °C; p_{Bc-SH}= 28.6 bar; - Pressure of Boiler C saturated steam: p
_{Bc-sat}= 33.6 bar; - Temperature and pressure of Boilers A and B: T
_{Bab}= 330 °C; p_{Bab}= 21.7 bar; - Boilers efficiency: η
_{Bab}= 0.58; η_{Bc}= 0.64; - Temperature after Boiler C tempering water addition: T
_{Bc-TEMP}= 370 °C; - Temperature loss from boilers A and B to 21.7 bar turbines: ∆T
_{Bab-turb}= 10 °C; - Isentropic efficiency of the turbines: η
_{IS-turb}= 0.60; - Isentropic efficiency of turbines for electrical power generation 4 and 5: η
_{IS-TG45}= 0.68; - Mechanical efficiency of the turbines (blades-shaft): η
_{B-S}= 0.98; - Electrical efficiency of the generators: η
_{el}= 0.95; - Turbine discharge pressure: p
_{VE}= 2.7 bar; - Nominal power of Turbo-fan of boilers: PVTIa = 253 kW; PVTIb = 201 kW; PVTIc = 615 kW;
- Nominal power of Turbo-generators: PTG1 = PTG2 = 1250 kW; PTG3 = 2500 kW; PTG4 = 3760 kW; PTG5 = 8510 kW.

_{el}, thermal efficiency η

_{th}, mechanical efficiency η

_{mech}, and global efficiency of the plant η

_{g}. By introducing the fuel heat power input of Boiler C (Q

_{FUEL-Bc}) and of Boilers A and B (Q

_{FUEL-Bab}), the steam mass rate produced by Boiler C (m

_{VAP-Bc}) and by Boilers A and B (m

_{VAP-Bab}), specific enthalpy difference of steam between outlet and inlet of Boiler C (∆h

_{BC}) and of Boiler A and B (∆h

_{Bab}), total heat fuel power input of the plant (Q

_{FUEL}), useful heat power used in the process of sugar and ethanol production (Q

_{US}), heat power used in juice heating (Q

_{J}), effect one evaporation (Q

_{EVAP-EFF1}), alcohol production (Q

_{PAC}), sugar drying (Q

_{D}), and the efficiency parameters were calculated as follows:

_{FUEL-Bc}= m

_{VAP-Bc}× ∆h

_{Bc}× η

_{Bc}(kW)

_{FUEL-Bab}= m

_{VAP-Bab}× ∆h

_{Bab}× η

_{Bab}(kW)

_{FUEL}= Q

_{FUEL-Bc}+ Q

_{FUEL-Bab}(kW)

_{US}= Q

_{J}+ Q

_{EVAP-EFF1}+ Q

_{PAC}+ Q

_{D}(kW)

_{mech}= P

_{mills}+ P

_{shredder}+ P

_{crusher}(kW)

_{g}= η

_{el}+ η

_{th}+ η

_{mech}(kW)

## 3. Results and Discussion

#### 3.1. Milling Rate Dependence

_{VAP-R28}, m

_{VAP-R21}). Under these conditions, the computed global efficiency of the plant is 58.49%. Notice the stream of 15.1 kg/s, and the one of 0.3 kg/s flowing through the 28.6 to 21.7 bar and 21.7 to 2.7 bar reduction stations, respectively. Figure 4 reports the steam mass rate and specific enthalpy in each step of the cogenerative cycle.

#### 3.2. Exergy Losses

_{II}) and the loss of energy related to critical steps. The η

_{II}in the steady-state current condition is equal to 19.62%. Equation 22 shows how to calculate η

_{II}, where P

_{el}stands for electrical power, P

_{mech}for mechanical power, Q for useful heat power, and Q

_{FUEL}is the fuel energy input.

_{REF}– T

_{REF}× (s – s

_{REF}) (kJ/kg)

_{LOSS}). Energy balance in a steady-state process is:

_{in}+ ε

_{th}= ε

_{out}+ P

_{mech}+ ε

_{LOSS}(kJ/kg)

#### 3.3. Renewable Efficiency

_{REN}) of the plant can be defined as the amount of energy produced in a renewable way with respect to the total energy required by the plant. Since bagasse represents 31% of the total cane weight, and its mean LHV (7984 kJ/kg) is known, it is possible to calculate the instantaneous bagasse flow (m

_{BAG}), the yearly bagasse production (m

_{BAG-y}), and the renewable efficiency, as shown in the following equations, by introducing the concepts of year equivalent hours HR

_{eq-y}, bagasse yearly available heat power Q

_{BAG-y}:

_{BAG}= 0.31× m

_{c}(kg/s)

_{BAG-y}= m

_{BAG-yx}× LHV

_{BAG}

_{REN}higher than 100%.

Process | Energy loss [%] |
---|---|

Boiler C tempering | 3.6 |

28.6 to 21.7 bar reduction | 3.2 |

21.7 to 2.7 bar reduction | 26.4 |

2.7 bar head tempering | 1.2 |

**Figure 6.**Global efficiency (as defined in Equation 21) of the sugarcane plant as a function of milling rate.

#### 3.4. Sensitivity Analysis

_{g}, and the selected variables were pressures and temperatures of boilers and tempering. The analysis was performed on the nominal plant operation. When a variable effect is studied all the others are kept constant. Figure 7a shows the effect of Boiler C outlet pressure on the η

_{g}. Notice that an optimum is detected.

_{g}decreases. The decrease in the pressure of Boilers A and B has the opposite effect on η

_{g}.

_{g}with the increase of inlet temperature of Boiler C with a fixed tempering temperature. The rise of the boiler internal temperature implies a mix of a high temperature steam with a lower temperature water. The bigger the temperature difference, the stronger the effect on η

_{g}, because more saturated steam has to be condensed, thereby losing important energy content. The temperature inside the boiler does not affect the steam reduction, because this variable only involves the thermal exchange inside the boiler, but the boiler’s output conditions are controlled by the temperature at tempering output, and are then fixed. Figure 8b shows the decreasing trend of global efficiency with the temperature of Boilers A and B. If the temperature raises, the enthalpy of steam increases, the change of operation point leads to an increase of enthalpy difference across the expansion. Thus, the steam mass rate required by the 21.7 bar turbines decreases and the steam mass rate through the 21.7 to 2.7 bar reduction station increases to fulfill the requirement of the process. If the temperature is reduced from the nominal one, no steam pressure reduction is necessary and the η

_{g}has a positive increasing trend with increasing temperature.

_{g}as displayed in Figure 9. Its range is limited by the boiler’s internal temperature. The tempering is aimed at protecting the turbines from thermal stress by controlling the temperature of steam, but it represents a considerable energy loss, as shown in Table 2. Losses decrease if the tempering temperature is closer to the boiler internal temperature.

#### 3.5. Numerical Optimization

_{g}was carried out with an algorithm based on the Pattern Search method [20,21].

_{g}(Equation 21), and the vector of variables was composed by pressures and temperatures of the boilers and temperature of Boiler C tempering (T

_{Bc.temp}). In practical cases, it has to be considered that all the machines and equipment are designed with specific nominal values of temperatures and pressures. In particular, it is not allowed to change pressures at the inlet of turbines. The tempering output temperature cannot be equal to the one inside the boiler because it would lose its function to prevent variations in boiler’s temperature to affect the turbines. For these reasons, it was chosen to have a minimum gap of 10 °C.

Temperature | Minimum value (°C) | Maximum value (°C) |
---|---|---|

Boiler C | 365 | 540 |

Boilers A and B | 320 | 340 |

Boiler C tempering output | 365 | 375 |

## 4. Scenarios to Improve the Global Efficiency

Milling rate: 430 Ton/h | p boiler C (bar) | T boiler C (°C) | p boilers A, B (bar) | T boilers A, B (°C) | T_{outlet} tempering boiler C (°C) | Global efficiency |
---|---|---|---|---|---|---|

Current case | 28.6 | 410 | 21.7 | 330 | 370 | 0.5849 |

Optimal case | 28.6 | 380 | 21.7 | 320 | 370 | 0.5880 |

#### 4.1. Option 1 Repowering

_{el-surplus}the total plant electrical power surplus, the energy index E

_{index}is expressed in Equation 31:

_{7}, has to be lower than the saturation temperature, and hence, a security margin of 3 °C was chosen. The steam fraction x

_{4}at the end of the expansion has to be higher than 0.87 to avoid excessive condensation inside the turbine, and the flow rates in the turbines have to be higher than 0. For each possible pressure and temperature combination of the boilers, pressure and mass rate of regeneration are optimized and verified to satisfy all the technical and thermodynamic constraints. The results are shown in Table 5. In Table 5, the minimum new boiler capacity and the power required by the new turbine are reported as well. These values are necessary for the design of the new equipment.

**Figure 11.**Effect of new cycle boiler’s pressure and temperature on energy surplus index in Option 1 repowering.

**(a)**Pressure;

**(b)**Temperature.

Boiler P (bar) | Boiler T (°C) | Boiler capacity (kg/s) | Boiler Inlet T (°C) | p reg (bar) | m_{vap} reg (kg/s) | P turbine top (MW) | η_{is} | P_{surplus} (MW) | η_{II} plant | Surplus (kWh/ton_{BAG}) |
---|---|---|---|---|---|---|---|---|---|---|

45.1 | 440 | 57.4 | 176 | 9.8 | 7.9 | 37.2 | 0.871 | 23.9 | 0.2371 | 179.2 |

64.7 | 485 | 56.7 | 187 | 12.5 | 8.6 | 39.9 | 0.863 | 26.5 | 0.2459 | 198.7 |

66.7 | 510 | 55.9 | 192 | 13.7 | 8.7 | 42.2 | 0.9000 | 28.7 | 0.2534 | 215.4 |

85.3 | 515 | 56.5 | 196 | 14.8 | 9.2 | 41.5 | 0.8468 | 27.9 | 0.2507 | 209.4 |

86.3 | 515 | 56.8 | 199 | 15.8 | 9.5 | 41.5 | 0.8440 | 27.8 | 0.2505 | 208.9 |

86.3 | 540 | 55.5 | 199 | 15.7 | 9.1 | 43.8 | 0.8858 | 30.2 | 0.2584 | 226.4 |

87.3 | 515 | 57.1 | 202 | 16.7 | 9.8 | 41.4 | 0.8411 | 27.8 | 0.2503 | 208.5 |

104 | 540 | 56.8 | 208 | 19.0 | 10.2 | 42.9 | 0.8415 | 29.2 | 0.2551 | 219.0 |

106.9 | 540 | 57.1 | 210 | 19.7 | 10.4 | 42.8 | 0.8350 | 29.0 | 0.2545 | 217.8 |

108.9 | 540 | 56.9 | 208 | 18.8 | 10.2 | 42.7 | 0.8306 | 28.9 | 0.2542 | 217.0 |

#### 4.2. Option 2 Repowering

#### 4.3. Economic Analysis of Repowering Options

Boiler P (bar) | Boiler T (°C) | Boiler capacity (kg/s) | Boiler Inlet T (°C) | p reheat (bar) | p reg (bar) | m_{vap} reg (kg/s) | P turbine top (MW) | η_{is} | P surplus (MW) | η_{II} plant | Surplus (kWh/ton_{BAG}) |
---|---|---|---|---|---|---|---|---|---|---|---|

45.1 | 440 | 51.9 | 163 | 22.2 | 6.3 | 5.7 | 39.2 | 0.9 | 25.9 | 0.2439 | 194.3 |

64.7 | 485 | 52.1 | 181 | 35.2 | 11.5 | 6.9 | 42.3 | 0.9 | 28.9 | 0.2540 | 216.8 |

66.7 | 510 | 49.8 | 176 | 30.9 | 8.4 | 6.1 | 43.3 | 0.9 | 29.9 | 0.2574 | 224.3 |

85.3 | 515 | 49.0 | 179 | 32.1 | 10.2 | 6.1 | 44.9 | 0.9 | 31.4 | 0.2625 | 235.6 |

86.3 | 515 | 50.9 | 190 | 40.5 | 11.8 | 7.3 | 45.1 | 0.9 | 31.6 | 0.2631 | 236.9 |

86.3 | 540 | 47.5 | 177 | 30.6 | 12.0 | 5.5 | 45.3 | 0.9 | 31.7 | 0.2637 | 238.1 |

87.3 | 515 | 52.3 | 221 | 32.9 | 22.1 | 9.3 | 45.0 | 0.9 | 31.5 | 0.2627 | 236.1 |

104.0 | 540 | 51.3 | 218 | 41.7 | 21.5 | 8.9 | 46.8 | 0.9 | 33.1 | 0.2683 | 248.5 |

106.9 | 540 | 52.6 | 226 | 47.8 | 24.3 | 9.8 | 46.9 | 0.9 | 33.2 | 0.2687 | 249.5 |

108.9 | 540 | 50.3 | 204 | 43.8 | 15.3 | 8.0 | 47.3 | 0.9 | 33.6 | 0.2699 | 251.9 |

_{w}of 600 L/s and global heat exchange coefficient U equal to 2550 W/m

^{2}·K are considered. The total purchased equipment of Option 1 and Option 2 cost 18.91 and 18.92 million Euro, respectively. Peters and Timmerhaus estimate that the cost of purchased equipment represents 15%–40% of total fixed capital cost [23]. Considering 27.5% as an average value, the investment for Options 1 and 2 are 68.78 and 68.87 million Euro, respectively.

Variable | Option 1 | Option 2 |
---|---|---|

Investment (€) | 18,914,782 | 18,940,182 |

Plant surplus energy index (kWh/ton_{BAG}) | 226.4 | 251.9 |

Electrical energy sell revenue (€/year) | 8,665,500 | 9,639,600 |

20 years NPV (€) | 54,376,553 | 68,128,494 |

PBT (year) | 10 | 9 |

IRR (%) | 11.0 | 12.7 |

## 5. Conclusion

## Nomenclature

## Symbols and acronyms:

BIGCC | Biomass Integrated Gasification Combined Cycle |

BPST | Backpressure Steam Turbine |

CEST | Condensing Extraction Steam Turbine |

E_{index} | Surplus energy index (kWh/tonBA) |

HR | Hours |

HRSG | Heat Recovery Steam Generator |

IRR | Internal Rate of Return (%) |

LHV | Low Heating Value (kJ/kg) |

LMTD | Log Mean Temperature Difference |

MR | Milling Rate (ton/day) |

NPV | Net Present Value (€) |

P | Power (kW) |

PAC | alcohol plant |

PBT | Payback time (years) |

PV | Present Value (€) |

Q | Heat power (kW) |

T | Temperature (°C) |

U | Global heat exchange coefficient (W/m ^{2}·K) |

Y | Electrical energy requirement (kWh/day) |

c | Cost of bagasse unit (€/ton) |

cap | Production capacity of the boiler (kg/s) |

h | Specific enthalpy (kJ/kg) |

i | Interest |

m | Mass rate (kg/s) |

mr | Milling rate (kg/s) |

p | Pressure (bar) |

q | Share of process heat |

s | Specific entropy (kJ/kg·K) |

saving | Money saving (€) |

tc | Tons of cane |

tfh | Tons of fiber per hour |

ε_{LOSS} | Loss specific exergy (kJ/kg) |

∆T | Temperature difference (°C) |

∆h | Enthalpy difference (kJ/kg) |

η | Efficiency |

## Subscripts:

Bab | Boilers A and B |

Bc | Boiler C |

Bc-sat | Boiler C saturated steam |

Bc-temp | Boiler C tempering water |

Bc-SH | Boiler C superheated steam |

Bab | Boiler A and B |

BAG | Bagasse |

BAG-y | Yearly bagasse production |

BLD | Blades |

B-S | Blade to shaft |

BOT | Lower pressure cycle |

C-y | Yearly milled cane |

D | Sugar drier |

EFF | Effect of the evaporator |

EVAP | Evaporation |

FUEL | Fuel input |

FUEL-Bc | Fuel input in Boiler C |

FUEL-Bab | Fuel input in Boiler A and B |

H2.7 | Head of 2.7 bar |

H21 | Head of 21.7 bar |

HP | High pressure |

II | Second law |

IS | Isentropic |

J | Juice heating |

LOSS | Loss |

LP | Low pressure |

MILL | Milling |

MP | Medium pressure |

OPT | Optimal |

PAC | Alcohol plant |

REF | Reference |

REN | Renewable |

SAT | Saturation |

SV | Saving of the variable to which it is referred |

TEMP | Tempering |

TG | Turbo-generator |

TOP | High pressure cycle |

TURB | Turbine |

TURB21 | Turbines 21.7–2.7 bar |

TW | Tempering water |

US | Useful |

VAP | Steam |

VAP-Bc | Produced by Boiler C |

VAP-Bab | Steam produced by Boiler A and B |

VAP-PAC | Steam demand of the alcohol plant |

VAP-N2.7 | Steam need at 2.7 bar |

VAP-N21 | Steam need at 21.7 bar |

VAP-N28 | Steam need at 28.6 bar |

VAP-R21 | Steam reduced from 21.7 to 2.7 bar |

VAP-R28 | Steam reduced from 28.6 to 21.7 bar |

VAP-reg | Steam for the regeneration |

VAP-TURB | Steam of turbines |

VAP-reg | Steam for the regeneration |

VAP-TURB | Steam of turbines |

VE | Turbine exhaust steam |

VTI | Turbo-fan |

W | Water |

el | Electrical |

eq-y | Year equivalent |

g | Global |

in | Inlet |

is | Isentropic |

mech | Mechanical |

mill | Mills |

need | requirement of the variable to which is referre |

out | Outlet |

reg | Regeneration |

shred | Shredder |

surplus | Surplus |

th | Thermal |

turb | Turbine |

vap | Steam |

vap-boiler | Capacity of the new boiler |

vapHP | High pressure steam |

vapMP | Middle pressure steam |

vapLP | Low pressure steam |

vapTOPprocess | Steam from high pressure cycle to the process |

## Author Contributions

## Conflicts of Interest

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## Share and Cite

**MDPI and ACS Style**

Colombo, G.; Ocampo-Duque, W.; Rinaldi, F.
Challenges in Bioenergy Production from Sugarcane Mills in Developing Countries: A Case Study. *Energies* **2014**, *7*, 5874-5898.
https://doi.org/10.3390/en7095874

**AMA Style**

Colombo G, Ocampo-Duque W, Rinaldi F.
Challenges in Bioenergy Production from Sugarcane Mills in Developing Countries: A Case Study. *Energies*. 2014; 7(9):5874-5898.
https://doi.org/10.3390/en7095874

**Chicago/Turabian Style**

Colombo, Guido, William Ocampo-Duque, and Fabio Rinaldi.
2014. "Challenges in Bioenergy Production from Sugarcane Mills in Developing Countries: A Case Study" *Energies* 7, no. 9: 5874-5898.
https://doi.org/10.3390/en7095874